Metal Machining - Theory and Applications Episode 1 Part 5 potx

The freshly formed, unoxidized, nature of the chip surface, created by the parting of the chip from the work typically less than 10–3s before it reaches the end of the contact length, an

and PE2in Figure 2.25 PE1represents theoretical analyses (Appendix 3) when the rough-ness is imagined to be on the tool surface and PE2when it is imagined to be on the chip

However, for large values of s/klocal, both regions have almost the same upper boundary,

with c (equation (2.26)) approximately equal to 1 One would then expect

s

klocal

In those circumstances, when m is measured to be < 1, this seems to be a reasonable

relation For example, in Figure 2.23, for the free machining steels when the rake face

temperature is below 600˚C, m is roughly the same as the ratio of m for the steel to that for

the plain carbon steel at the same temperature However, equation (2.28) cannot explain

observations of m > 1, of the sort recorded in Figure 2.23(b) for the non-free machining

steel or for the free machining steels above 600˚C

Friction coefficients greater than 1.0

The plastic contact mechanics modelling reviewed in Appendix 3, which leads to c≤ 1, for the most part assumes that the asperity does not work harden and that the load on the asper-ity is constant through its make and break life cycle In the final section of Appendix 3 there is a brief speculation about departures from these assumptions that could lead to

larger values of c and to m > 1 All proposals require the shear strength of the junction to

be maintained while the normal stress is unloaded It is certain that, for this to occur, the strongest levels of adhesion must exist between the asperities and the tool The freshly formed, unoxidized, nature of the chip surface, created by the parting of the chip from the work typically less than 10–3s before it reaches the end of the contact length, and the high temperatures reached at high cutting speeds, are just the conditions that could promote strong adhesion (or friction welding) However, there is, at the moment, no quantitative theory to relate friction coefficients greater than 1 to the underlying asperity plastic prop-erties and state of the interface

The proper modelling of friction is crucial to the successful simulation of the machin-ing process This section, with Appendix 3, is important in settmachin-ing current knowledge in a contact mechanics framework, but there is still work to be done before friction in metal machining is fully understood

2.4.2 Lubrication in metal cutting

The previous section has emphasized the high friction conditions that exist between a chip and tool, in the absence of solid lubricants The conditions that lead to high friction are

Friction, lubrication and wear 73

Table 2.4 Tool surface roughness and contact stress severity data

10klocal/E*,°

high cutting speeds – for steels, speeds greater than around 100 m/min when the feed rate

is 0.1 to 0.2 mm However, earlier in this chapter (Figure 2.7) liquid lubrication was demonstrated at low cutting speeds; and one of the earliest questions asked of metal cutting (Section 2.1) was how can lubricant penetrate the rake face contact?

The question can now be asked in the context of the contact mechanics of the previ-ous section Figure 2.27 shows, somewhat schematically, the contact between the chip and tool The hatched region represents the real area of contact, covering 100% of the contact near the cutting edge, where the normal stress is high, and reducing to zero towards the end of the contact It is now generally agreed that neither gaseous nor liquid lubricants can penetrate the 100% real contact region, but they can infiltrate along the non-contact channels at the rear of the contact These channels may typically be from half

to one chip thickness long, depending on the normal contact stress distribution (Figure 2.22) Their height depends on the surface roughness of the cutting tool, but is typically

0.5 to 1 mm (Figure 2.26) If the lubricant reacts with the chip to reduce friction in the

region of the channels, the resistance to chip flow is reduced, the primary shear plane angle increases, the chip becomes thinner and unpeels from the tool Thus, a lubricant does not have to penetrate the whole contact: by attacking at the edge, it can reduce the

whole So the question becomes: what is the distance lp(Figure 2.27) that a gas or liquid can penetrate along the channels? The following answer, for the penetration of gaseous

oxygen and liquid carbon tetrachloride along channels of height h, is based on work by

Williams (1977)

It is imagined that the maximum penetration results from a balance of two opposing transport mechanisms: the motion of the chip carrying the gas or liquid out of the contact and the pressures driving them in For a gas, absorption on to the back of the freshly formed chip is the mechanism of removal from the contact The absorption creates a gas pressure gradient along the channel which drives the gas in Williams identified two mechanisms of inward flow, based on the kinetic theory of gases: viscous (Poiseuille) flow at high gas vapour pressure and Knudsen flow at low pressures, when the mean free

path of the gas is greater than the channel height h He showed that l (mm) is inversely

74 Chip formation fundamentals

Fig 2.27 (a) Defining the penetration distance lpof the lubricant into the rear of the contact region and (b) derived feed/speed regions of complete and negligible penetration, for oxygen

proportional to the chip velocity Uchip(m/min) with the constant of proportionality

depending on the gas molecular diameter, obtained from its molecular weight M and its density in the liquid state rliquid(kg/m3), on its vapour pressure pv(Pa), its viscosity h (Pa s) absolute temperature qTand on the height h (mm) For a channel much wider than its

height

h3 p2v M 2/3

lpUchip = the larger of 3.3 × 10–10— —— (——— ) (Poiseuille) (2.29a)

h qT rliquid

or

q3

T r4 liquid

For oxygen, at its normal partial pressure in air of ≈ 2 × 104Pa, and M = 32, rliquid= 1145 kg/m3, h = 20 × 10–6Pa s, qT= 293 and for h = 0.5 mm,

This is about half the value given by Williams, because of different assumptions about the cross-sectional shape of the channels; and it does depend strongly on the assumed value of

h.

Because of volume conservation, the product of Uchipand chip thickness t is the same

as of Uworkand feed f Equation (2.30a) can therefore be modified to

lp

t

At feeds and speeds for which lp/t is calculated to be > 1, total penetration of oxygen into the channels is expected When lp/t < 0.1, penetration may be considered negligible.

Figure 2.27 marks these regions as possibly lubricated, and not lubricated, respectively

It is important because it shows a size effect for the effectiveness of lubrication Williams (1977) also considered the penetration of liquids into the contact, driven by capillary forces and retarded by shear flow between the chip and the tool For carbon tetrachloride liquid (which also has a significant vapour phase contribution to its penetration) he concluded the limiting feeds and speeds for lubrication were about the same as for oxygen

Although it is certain that there can be no lubrication in the ‘no lubrication’ region of Figure 2.27, it is not certain that there will be lubrication in the ‘possible lubrication’ region Whatever penetrates the channels must also have time to react and form a low

fric-tion layer The time to react has also been studied by Williams (Wakabayashi et al., 1995).

It seems that this, rather than the ability to penetrate the channels, can be the controlling step for effective lubrication

It is not the purpose of this section to expand on the effectiveness of different lubricat-ing fluids for low speed applications This has been covered elsewhere, for example Shaw (1984) Rather, it is to gain an understanding of the inability of liquids or gases to influ-ence the contact at high cutting speeds The reason why cutting fluids are used at high speeds is to cool the work material and to flush away swarf

Friction, lubrication and wear 75

2.4.3 Wear in metal cutting

Finally, the sliding of the chip over the rake face, and of the work past the flank, causes the tool to wear away Tool wear will be considered in detail in Chapter 4 Here, the purpose is briefly to review knowledge of wear from other studies, to create a standard to which tool wear can be related

One of the most simple types of wear test is a pin on disc test (Figure 2.28) A

cylin-drical pin of cross-section A is pressed with a load W against a rotating disc which has some sliding speed U against the pin The rate of loss of height, h, of the pin is measured

against time Usually there is an initial, running-in, time of high wear rate, before a constant, lower, rate is established A common observation is that, in the steady state, the

wear volume rate, Adh/dt in this example, is proportional to W and the sliding speed.

Archard’s wear law (Archard and Hirst, 1956) may be written

where the constant of proportionality kswris called the specific wear rate and has units of

inverse pressure (In the wear literature kswris written k, but k has already been used in this

book for a metal’s shear flow stress.)

The proportionality of wear rate to load and speed is perhaps obvious However,

Archard considered the mechanics of contact to establish likely values for kswr He consid-ered two types of contact, abrasive and adhesive (Figure 2.29) – the terminology is expanded on in Appendix 3 In the abrasive case, the disc surface consists of hard, sharp conical asperities (as might be found on abrasive papers or a grinding wheel) They dig

76 Chip formation fundamentals

Fig 2.28 A pin on disc wear test and a typical variation of pin height with time

into the softer pin to create a number of individual real contacts, each of width 2rr As a

result of sliding, a scratch is formed of depth rrtanb, where b is the slope of the cones If

it is supposed that all the scratch volume becomes wear debris, the volume wear per unit

time is Ur2

rtan b At the time Archard was writing, the analogy was made between the

indentation of the cone into the flat and a hardness test, to relate the contact width to the

load W on the cone Noting that, during sliding, the load W is supported on the semicircle

of area pr2

r/2, r2

r was equated to (2/p)(W/H), where H is approximately the Vickers or

Brinell hardness of the softer surface By substituting this into the expression for the scratch volume and summing over the large number of scratches that contribute to the wear process, it is easy to convert equation (2.31a) to the form of (2.31b), where a

dimension-less wear coefficient K has been introduced instead of the specific wear rate kswr, with a magnitude as written for this abrasive example

A similarly simple model for adhesive wear (also Figure 2.29) assumes that a

hemi-spherical wear particle of radius rris torn from the surface every time an asperity slides a

distance 2rr, and that the real contact pressure is also H It leads to the adhesive wear esti-mate of K also being included in equation (2.31b)

— = — snU; K = ——— for abrasive wear

(2.31b) 1

= — for adhesive wear 3

If these equations were being derived today, there would be discussion as to whether

the real contact pressure was H (equivalent to 5k) or only to k (Section 2.4.1 and Appendix 3) However, such discussion is pointless It is found that the K values so deduced are

orders of magnitude different from those measured in experiments Actual wear mecha-nisms are not nearly as severe as imagined in these examples Different asperity failure mechanisms are observed, depending on the surface roughness, through the plasticity

index already introduced in Section 2.4.1 and on the level of adhesion expressed as s/k or

m Figure 2.30 is a wear mechanism map showing what failure mode occurs in what

condi-tions It also shows what ranges of K are typical of those modes (developed from Childs,

1980b, 1988)

Friction, lubrication and wear 77

Fig 2.29 Schematic views of abrasive and adhesive wear mechanisms

The initial wear region is the running-in regime of Figure 2.28 Surface smoothing occurs until the contacting asperities deform mainly elastically If the surface adhesion is small

(mild wear region), material is first oxidized before it is removed – values of K from 10–4to

10–10are measured (all the data are for experiments in air, nominally at room temperature)

At higher adhesions subsurface fatigue (delamination) is found, with K around 10–4 Sometimes, running-in does not occur and surfaces do tear themselves apart (severe adhesive

wear), but even then K is found to be only 10–2to 10–3, compared with the value of 1/3

predicted above Finally, if abrasive conditions do exist, K is found between 10–1and 10–4, depending on whether the abrasive is fixed on one surface (2-body) or is loose (3-body) What is the relevance of this to metal machining? In Chapter 1, it was described how the economics of machining lead to the use of, for example, cemented carbide tools at cutting speeds and feeds such that the tools last only 5 to 10 minutes before wearing out Definitions of wear-out differ from application to application, but common ones are that the

flank wear length is less than 300 mm, or that the depth of any crater on the rake face is less than 60 mm Figure 2.31(a) shows a worn tool, with crater depth hcand flank depth wear hf

hfis related to the length of the wear land by tan g, where g is the flank clearance angle.

Figures 2.31(b) and (c) are examples of wear measured for a low alloy steel at a feed of 0.12

mm and a cutting speed of 225 m/min, which is near the economic speed For the flank,

dhf/dt ≈ 2 mm/min; for the crater example dhc/dt ≈ 7 mm/min Supposing the contact stress level is characterized by sn/k ≈ 1, and noting that H ≡ 5k, values of K, from equation

2.31(b), are 4 × 10–8on the flank, up to 3 × 10–7on the rake (the speed of the chip was

78 Chip formation fundamentals

Fig 2.30 A wear mechanism map

half that of the work) Considering that s/k is large in machining, these values are smaller

than expected from the general wear testing experience summarized in Figure 2.29 (There

is another point: the proportionality between dh/dt and sn/k in equation (2.31) is only established for conditions in which Ar/An< 0.5 Values larger than this occur over much

of the tool contacts in machining However, the uncertainty that this places in the deduced

values of K is not likely to alter the orders of magnitude deduced for its values.)

There is one point to be made: the K values in Figure 2.30 are appropriate for the wear

of the chip and work by the tool, rather than of the tool by the chip or work! In Figure 2.30, the plasticity index is, in effect, the ratio of the work material’s real contact stress to its shear flow stress To use the map to determine wear mechanisms in the tool, it seems appropriate to redefine the index as the ratio of the contact stress in the work to the tool

material’s shear flow stress For typical tool materials (HV = 10 GPa to 15 GPa) and work materials (say HV = 2.5 GPa), this would effectively reduce the plasticity index value for

the tool about fivefold relative to the work For typical work plasticity index values of about 20 (Table 2.4), this would place the tool value at about 4, in the elastic range of Figure 2.30 The mechanisms available for tool wear are likely to be fatigue and chemical reaction (oxidation) with the atmosphere

This conclusion is based on a continuum view of contact mechanics In practice, work materials contain hard abrasive phases and tool materials contain relatively soft binding phases, so abrasion occurs on a microstructural scale The transfer of work material to the tool, by severe adhesive wear, can also increase the tool stresses At the temperature of cutting, chemical reactions can occur between the tool and work material as well as with the atmosphere The story of abrasive, mechanical fatigue, adhesive and reaction wear of cutting tools is developed in Chapter 4

2.5 Summary

The sections of this chapter have established the severe mechanical and thermal conditions typical of machining A certain amount of factual information has been gathered and deductions made from it, but for the most part this has been at the level of observation Predictive mechanics is taken up in the second half of this book, from Chapters 6 onwards

Summary 79

Fig 2.31 (a) Flank and crater tool wear regions, with typical (b) flank and (c) crater wear observations

First however, materials aspects of, and experimental techniques for, machining studies are introduced in Chapters 3 to 5

References

Archard, J F and Hirst, W (1956) The wear of metals under unlubricated conditions Proc Roy Soc.

Lond A236, 397–410.

Boothroyd, G and Knight, W A (1989) Fundamentals of Machining and Machine Tools New York:

Marcel Dekker.

Boston, O W (1926) A research in the elements of metal cutting Trans ASME 48, 749–848.

Chandrasekeran, H and Kapoor, D V (1965) Photoelastic analysis of tool-chip interface stresses.

Trans ASME J Eng Ind 87B, 495–502.

Childs, T H C (1972) The rake face action of cutting lubricants Proc I Mech E Lond 186,

717–727.

Childs, T H C (1980a) Elastic effects in metal cutting chip formation Int J Mech Sci 22,

457–466.

Childs, T H C (1980b) The sliding wear mechanisms of metals, mainly steels Tribology

International 13, 285–293

Childs, T H C (1988) The mapping of metallic sliding wear Proc I Mech E Lond 202 Pt C,

379–395.

Childs, T H C and Maekawa, K (1990) Computer aided simulation of chip flow and tool wear.

Wear 139, 235–250.

Childs, T H C., Richings, D and Wilcox, A B (1972) Metal cutting: mechanics, surface physics

and metallurgy Int J Mech Sci 14, 359–375.

Eggleston, D M., Herzog, R and Thomsen, E G Some additional studies of the angle relationships

in metal cutting Trans ASME J Eng Ind 81B, 263–279.

Herbert, E G (1928) Report on machinability Proc I Mech E London ii, 775–825.

Kato, S., Yamaguchi, Y and Yamada, M (1972) Stress distribution at the interface between chip and

tool in machining Trans ASME J Eng Ind 94B, 683–689.

Kobayashi, S and Thomsen, E G (1959) Some observations on the shearing process in metal

cutting Trans ASME J Eng Ind 81B, 251–262.

Lee, E H and Shaffer, B W (1951) The theory of plasticity applied to a problem of machining.

Trans ASME J Appl Mech 18, 405–413.

Mallock, A (1881–82) The action of cutting tools Proc Roy Soc Lond 33, 127–139.

Merchant, M E (1945) Mechanics of the metal cutting process J Appl Phys 16, 318–324.

Oxley, P L B (1989) Mechanics of Machining Chichester: Ellis Horwood.

Shaw, M C (1984) Metal Cutting Principles, Ch 13 Oxford: Clarendon Press.

Shirakashi T and Usui, E (1973) Friction characteristics on tool face in metal machining J JSPE

39, 966–972.

Taylor, F W (1907) On the art of cutting metals Trans ASME 28, 31–350.

Trent, E M (1991) Metal Cutting, 3rd edn., Ch.9 Oxford: Butterworth Heinemann.

Tresca, H (1878) On further applications of the flow of solids Proc I Mech E Lond pp 301–345

and plates 35–47.

Wakabayashi, T., Williams, J A and Hutchings I M (1995) The kinetics of gas phase lubrication in

the orthogonal machining of an aluminium alloy Proc I Mech E Lond 209Pt.J, 131–136 Weiner, J H (1955) Shear plane temperature distribution in orthogonal cutting Trans ASME 77,

1331–1341.

Williams, J A (1977) The action of lubricants in metal cutting J Mech Eng Sci 19, 202–212.

Zorev, N N (1966) Metal Cutting Mechanics Oxford: Pergamon Press.

80 Chip formation fundamentals

Work and tool materials

In Chapter 2, the emphasis is on the mechanical, thermal and friction conditions of chip formation The different work and tool materials of interest are introduced only as exam-ples In this chapter, the materials become the main interest Table 3.1 summarizes some

of the main applications of machining, by industrial sector and work material group, while Table 3.2 gives an overview of the classes of tool materials that are used In Section 3.1 data will be presented of typical specific forces, tool stresses and temperatures generated when machining the various work groups listed in Table 3.1 In Section 3.2 the properties

of the tools that resist those stresses and temperatures will be described

A metal’s machinability is its ease of achieving a required production of machined components relative to the cost It has many aspects, such as energy (or power) consumption, chip form, surface integrity and finish, and tool life Low energy consumption, short (broken) chips, smooth finish and long tool life are usually aspects of good machinability Some of these aspects are directly related to the continuum mechanical and thermal conditions of the

Table 3.1 Some machining activities by work material alloy and industrial sector

Carbon and Structures Power train, Power train, Structures Printer

hydraulics hydraulics gear

fasteners

Aluminium Structures Engine block Airframe For corrosion Scanning

and pistons spars, skins resistance mirrors, disc

substrates

resistance

blades and exchangers, discs and corrosion

resistance

airframe resistance

machining process In principle, they may be predicted by mechanical and thermal analy-sis (but at the current time some are beyond prediction) Other aspects, principally tool life, depend not only on the continuum surface stresses and temperatures that are generated but also on microstructural, mechanical and chemical interactions between the chip and the tool Table 3.3 summarizes these relations and the principal disciplines by which they may

be studied (perhaps chip/tool friction laws should come under both the applied mechanics and materials engineering headings?) This chapter is mainly concerned with the work material’s mechanical and thermal properties, and tool thermal and failure properties, which affect machinability Tool wear and life are so important that a separate chapter, Chapter 4, is devoted to these subjects

3.1 Work material characteristics in machining

According to the analysis in Chapter 2, cutting and thrust forces per unit feed and depth of cut, and tool stresses, are expected to increase in proportion to the shear stress on the

primary shear plane, other things being equal This was sometimes written k and some-times kmax

Forces also increase the smaller is the shear plane angle and hence the larger is the strain in the chip The shear plane angle, however, reduces the larger is the strain

harden-ing in the primary shear region, measured by Dk/kmax (equation (2.7)) Thus, kmax and

Dk/kmaxare likely to be indicators of a material’s machinability, at least as far as tool forces and stresses and power consumption are concerned Figure 3.1 gathers information on the typical values of these quantities for six different groups of work materials that are impor-tant in machining practice The data for steels exclude quench hardened materials as, until

82 Work and tool materials

Table 3.2 Recommended tool and work material combinations

√ good; O all right in some conditions; ⊗ possible but not advisable; x to be avoided.

Table 3.3 Mechanical, thermal and materials factors affecting machinability

Main tools for study Process variables Machinability attribute

Applied mechanics Work mechanical and thermal properties Power consumption

and thermal analysis Tool thermal properties Tool stresses and temperatures

Tool failure properties Tool failure Chip/tool friction laws Surface integrity and finish Materials engineering Work/tool wear interactions Tool wear and life